X65 bimetallic sheet

Accepted Manuscript Study on the microstructure and mechanical properties of explosive welded 2205/X65 bimetallic sheet Lin-Jie Zhang, Qiang Pei, Jian-Xun Zhang, Zong-Yue Bi, Ping-Cang Li PII: DOI: Reference:

S0261-3069(14)00629-3 http://dx.doi.org/10.1016/j.matdes.2014.08.013 JMAD 6710

To appear in:

Materials and Design

Received Date: Accepted Date:

21 June 2014 6 August 2014

Please cite this article as: Zhang, L-J., Pei, Q., Zhang, J-X., Bi, Z-Y., Li, P-C., Study on the microstructure and mechanical properties of explosive welded 2205/X65 bimetallic sheet, Materials and Design (2014), doi: http:// dx.doi.org/10.1016/j.matdes.2014.08.013

This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.

Study on the microstructure and mechanical properties of explosive welded 2205/X65 bimetallic sheet Lin-Jie Zhang a , Qiang Pei a, Jian-Xun Zhang a, Zong-Yue Bi b, Ping-Cang Li c a

State NH\ODERUDWRU\RIPHFKDQLFDOEHKDYLRUIRUPDWHULDOV;L¶DQ-LDRWRQJ8QLYHUVLW\;L¶DQ&KLQD

b

Baoji Petroleum Steel Pipe Co., Ltd., Baoji, Shaanxi, 721008, China

c

;L¶DQ Tianli Clad Metal Materials Co., Ltd.;L¶DQ710201, China

Abstract: In this study, microstructural inhomogeneity and mechanical properties of explosive welded 2205 stainless steel/X65 pipe steel bimetallic sheets were investigated. The explosion-bonded 2205/X65 bimetallic sheets had good shear strength. The tensile shear fracture primarily occurred in the interior of X65 material and primarily exhibited dimple morphology. Fine crystal grains in the 0.5-2 micron range were found all over the narrow localized melted zone near the 2205/X65 interface, whereas a coarse columnar crystal structure growing along the perpendicular direction to the interface formed in the wider localized melted zone. Quasi-cleavage fracture morphology was observed in the coarse columnar crystal region after the stratified tensile test. The junction of the three regions near the interface with large differences in morphology was the weak point in the bimetallic sheet, where Y-shaped cracking easily occurred under a loading force. Stratified tensile test and micro-hardness tests for the explosively welded bimetallic sheet showed that severe hardening occurred in the 2205 cladding, and the most severe metal hardening occurred near the interface. Tests for 45° face bending and root bending tests were conducted under extreme conditions. The results showed that voids were prone to appeared in the peninsula and island morphologies near the interface.

Keywords: Bimetallic sheet; Explosive welding; Microstructure; Mechanical property; Inhomogeneity

1. Introduction Bimetallic sheets can fully exploit the respective advantages of the two materials that are used for a flyer plate and a parent plate, thereby achieving a performance that a single metal cannot provide. Thus, low cost can be seamlessly combined with high performance, offering a wide range of prospects for development. Explosive welding is one of the most widely used methods for manufacturing bimetallic sheets. Explosive welding can form a wavy metallurgical bonding interface between two metals without causing significant crystallization or phase transition.

Corresponding author. Tel.: 86-29-82663115; Fax: 86-29-82663115 E-mail: [email protected] (L-J Zhang)

1

Currently, researchers have successfully achieved the explosive welding of hundreds of combinations of similar or dissimilar metals. Miao et al. showed that medium steel/mild steel and stainless steel/mild steel bimetallic plates could be produced simultaneously by using honeycomb structure explosives and double sided explosive cladding [1]. Guo et al. reported that aluminium/316L stainless steel bimetallic pipe with excellent metallurgical bonding quality could be achieved by explosive cladding. They claimed that the aluminium/316L bimetallic pipe produce by explosive welding could endure the second plastic forming [2]. Zhou et al. produced the steel-fiber reinforced steel/aluminum bimetallic plates and discussed the effects of layer thickness distribution and fibre density on the ballistic resistance of steel fiber reinforced steel/aluminum bimetallic plates [3]. Raghukandan explosively cladded low carbon steel with copper and discussed the effects of flyer thickness, loading ratio, angle of inclination and stand-off on the clad strength [4]. Ahmet et al. experimentally demonstrated that copper and stainless steel could be bonded with a good quality of bonding properties through explosion welding [5]. Bina et al. produced copper/austenitic stainless steel couple by explosive welding and improved its mechanical properties by post weld heat treatment [6]. Gülenc showed that aluminium could be well bonded to copper sheet by using explosive welding [7]. In Kahraman et al.¶s study, Ti6Al4V alloy plates and commercial copper plates were bonded through explosive welding process [8]. Xia et al. investigated the microstructure and mechanical properties of TA2/2A12 bimetals fabricated by explosive welding [9]. In Bataev et al.¶s work, 10 plates of cp-titanium and 11 plates of cp-aluminum were explosively welded together to fabricated a 21-layer metallic-intermetallic laminate composites. The results showed that the combination of explosion welding and subsequent annealing was an effective and relatively inexpensive technology of metallic-intermetallic laminate composites fabrication [10]. Nizamettin et al. produced Ti6Al4V/304 stainless steel bimetallic plates through explosive welding process. No joining fault was formed in Ti6Al4V/304 interface, and also no melting voids and intermetallic compounds were observed [11]. Song et al. studied the microstructure of explosive cladding joints formed among parallel Ti and steel plates [12]. Wang et al. found that the tube of titanium and the bar of NiCr alloy could be bonded through explosive cladding technique with a good quality bonding [13]. Manikandan studied the effect of employing an interlayer of thin stainless steel in explosive welding of pure titanium/stainless steel bimetallic plates on the formation of intermetallic layer at the interface and the dynamic strength of the joint [14]. In Liu et al.¶s work, a defect-free metallurgical bonding between the Fe-based metallic glass foil and the aluminum plate was successfully obtained by explosive welding technique [15]. Zareie Rajani HR et al. compared the effect of fusion cladding and explosive cladding procedures on corrosion behavior of Inconel 625 cladding on plain carbon steel as substrate and found that corrosion resistance of fusion cladding Inconel 625 was weaker than explosive cladding superalloy [16]. Gerland et al. proved that a 2

primary explosive could properly clad a thin nickel film with thicknesses varying from 50 to ȝP to a thick aluminum alloy substrate with a good quality bonding [17]. Gong et al. explosively welded Cu-Al-Mn and QBe2 alloy and evaluated the microstructure and properties of the composite laminate obtained [18]. The increasing demand for oil and natural gas has, in turn, created an increasing demand for the long distance transportation and refining of highly corrosive crude oil and natural gas. Bimetallic sheets in which stainless steel is used as the material of the flyer plate instead of standard vessel steel can significantly improve the life of the pipeline and refining equipment. In recent years, scholars have conducted many studies on the microstructure, mechanical properties and failure mechanisms of stainless steel/carbon steel bimetallic sheets. Kacar et al. prepared bimetallic sheets by the explosive welding of 316L austenitic stainless steel and DIN-P355GH vessel steel and investigated the microstructure, hardness, tensile shear strength, and fracture toughness of the obtained bimetallic sheet [19]. The impact toughness of the bimetallic sheet was found to be higher than that of DIN-P355GH vessel steel. The authors submitted that the local melting region near the interface dominated the impact on the tensile shear strength of the interface. Kaya et al. studied the impact of the explosive ratio on the microstructure and the mechanical properties of a bimetallic sheet obtained by explosive welding of 316L austenitic stainless/Grade A ship steel [20]. A microstructure examination showed that increasing the explosive ratio transformed the cladding interface from a smooth to a wavy shape. Increasing the explosive ratio also increased the wavelength and amplitude of the waviness. The grains near the interface were elongated along the explosion direction. The hardness values of the explosively cladded joint increased with the explosive ratio. No separation was observed in the joint interface of the explosively cladded sheet after a three-point bending test. Mendes et al. reported on the effect of explosive characteristics on the welded interface between stainless steel AISI 304L and low alloy steel 51CrV4 in a cylindrical configuration [21]. The authors showed that the type of explosion and the type and proportion of the explosive sensitizers affected the main welding parameters, particularly the collision point velocity. All welded interfaces that were produced using an ammonium nitrate-based emulsion exhibited a localized melted and solidified zone with a chemical composition that depended on both the flyer and base metals. These melted zones could become incipient or even be eliminated using an ammonium nitrate fuel oil explosive. Zamani et al. studied the explosive welding process for bimetallic corrosion resistant steel pipes in which the outer and inner pipes were CK22 carbon steel and 316L stainless steel, respectively [22]. In this study, the impact velocity of the pipes was the most important collision parameter and was calculated using a finite element simulation. A weldability window was developed for the stainless steel pipe (the inner pipe) and the carbon steel pipe (the outer pipe). To investigate the effects of the explosive welding process variables on the strength of the bond, Akbari Mousavi SAA et al. carried out a series of 3

explosive welding trials [23]. There was very little difference between the shear strength parallel and perpendicular to the detonation direction. In general, the bonding strength increased slightly with the stand-off distance. The results of a peel test showed that the interface peel toughness decreased when melting was present in the interface. A thicker melted interface corresponded to lower interfacial peel toughness. The shear test results showed that the shear strength decreased slightly when melted liquid was present at the interface. The results obtained in this study showed that the minimum and maximum interfacial fracture toughness values were obtained for a vortex-shaped interface and a straight interface, respectively. The authors emphasized that the peel toughness could not be predicted using the traditional ASTM shear test. Akbari Mousavi SAA et al. numerically analyzed the strain levels induced in the plates and the direction of the shear stress in the collision zone in the explosive welding process [24]. In a study of Acarer et al., different welding interfaces (i.e., straight, wavy and continuous solidified-melted interfaces) were obtained by changing the explosive welding parameters such as the stand-off distance, the explosive loading and the anvils [25, 26]. The tensile shear and bending test results showed that heat-treated specimens had more strength than specimens that were not subjected to heat treatment. The tensile shear test results showed that flat and wavy interfaces provided approximately the same shear strength. The bending zone exhibited some cracks after bending tests were conducted on the unheated specimens. X65 pipe steel offers high strength, high toughness and a minor Bauschinger effect, resulting from thermo-mechanical controlled rolling. Using pipe steels such as X65 in the production of long-distance, high-pressure oil or gas transportation pipelines can reduce the overall project cost because smaller wall thicknesses can be used. The yield strength of 2205 duplex stainless steel is approximately twice that of standard austenitic stainless steel. The pitting resistance, stress corrosion and corrosion fatigue performance of 2205 duplex stainless steel are significantly better than those of conventional austenitic stainless steel. The linear expansion coefficient of 2205 duplex stainless steel is lower than that of austenitic stainless steel and close to that of carbon steel. Therefore, 2205 duplex stainless steel is particularly suitable for the production of the stainless steel/steel bimetallic sheet for use in oil and gas transportation pipes and refining equipment. Kacar et al. showed that the mechanical and corrosion properties of DIN-P355GH grade vessel steel can be increased by explosive cladding with 2205 grade duplex stainless steel [27]. The tensile shear strength test results showed that the duplex stainless steel-vessel steel cladded materials exhibited acceptable joint strengths. The hardness of the base and flyer plates near the bonding interface were similar, whereas the hardness of both the base and flyer plates generally increased at approximately ȝPDZD\IURPthe interface. The impact toughness of the cladded materials at a given test temperature was found to be significantly higher than that of the base material alone because of the high impact toughness of the duplex stainless steel layer. 4

In this study, large-scale 2205/X65 bimetallic sheet was fabricated by the explosive welding technique and its microstructure and mechanical properties were studied through optical microscopy, scanning electron microscopy (SEM), energy spectrum analysis (EDS), micro-hardness test, tensile test, tensile shear test and bending test. Compared with the previous studies about explosive welded bimetallic materials, this work focused more on the inhomogeneity of both microstructure and mechanical properties of the obtained 2205/X65 bimetallic sheet. The 2205/X65 bimetallic sheets discussed in this work would be used to manufacture long oil and gas pipelines and therefore would subjected to multiple bending process (i.e. JCOE process). The deformation and strain rate of bimetallic sheet would be highly heterogeneous during JCOE forming process and this might be beneficial to local shear failure. In order to minimize the risk of local failure, optimization of JCOE processing parameters based on numerical study was needed. Therefore, one of the major goals of this study is to provide detailed information about the mechanical properties of the bimetallic sheet for establishing a more realistic and practical numerical model of JCOE processing. In addition, the weak positions near the interface of the 2205/X65 bimetallic sheet were revealed through tensile tests and bending tests under extreme conditions, which provided the clues in optimizing the welding parameters of 2205/X65 bimetallic sheets. 2. Materials and methods 2.1 Materials and explosive welding process The flyer and base plates were made of 2205 duplex stainless steel and X65 pipe steel, respectively. The chemical compositions of the flyer and base plates are given in Table 1. The dimensions of the flyer and base plates were 6000 mm×2000 mm×2 mm and 6000 mm×2000 mm×16 mm, respectively. The explosive material was an ammonium nitrate fuel oil mixture. A multi-section arrangement of the explosive materials was used to control the spatial variation in the welding quality. The final thickness of the explosively bonded 2205/X65 bimetallic sheets was approximately 18 mm. Ultrasonic inspections were carried out for the explosively bonded 2205/X65 bimetallic sheets, and the results showed that the welding quality met the A grade requirements of ASTM: B898-11. 2.2 Metallographic studies The cross-section of the 2205/X65 bimetallic sheet was ground, polished and etched. The etchant consisted of 100 mL alcohol, 100 mL HCl and 5 g of CuCl2. A Nikon Eclipse MA200 type optical microscope was used for the microscopic studies. Higher resolution examination of the samples was carried out using an LS-JLLH-22 scanning electron microscope. Energy-dispersive X-ray spectrometry (EDS) analysis was used to characterize the distribution of the alloy elements across the 2205/X65 interface. 5

2.3 Mechanical tests The micro-hardness test was carried out using a load of 300 gf and a holding time of 10 s. The specimens for the tensile shear test were prepared parallel to the explosion direction from the 2205/X65 bimetallic sheet according to GB/T6396-2008 [28]. The tensile shear test was conducted on a universal mechanical testing machine for a shearing speed of 0.2 mm/min. Six slices of metallic materials with limited thicknesses of approximately 1.2 mm were cut from the explosively bonded 2205/X65 sheet in the horizontal direction, and a tensile test was performed on each slice of metal. Three face bending specimens with 8-mm thicknesses were tested, and the 2205/X65 interface in the face-bended specimens was examined. Finally, both face and root bending of the full thickness specimen were carried out under extreme conditions to reveal the weak points in the 2205/X65 bimetallic sheet. 2.4 Fracture observation The fractures obtained from both stratified tensile tests and tensile shear tests were examined using a LS-JLLH-22 scanning electron microscope. The fracture characteristics of the 2205/X65 bimetallic sheet were analyzed by combining the EDS test results for the alloying element contents on the fracture surface. 3. Results and discussion 3.1 Optical microstructure Metallographic studies showed that the metallurgical bonding of 2205 duplex stainless steel and X65 pipe steel was achieved by explosive welding, and the bonding interface of 2205/X65 bimetallic sheet had a wavy morphology, as shown in Fig. 1. It was reported that with the increasing of explosive loading the welding interface changed from straight to wavy interface [25, 26]. The morphology of the interface given in Fig. 1 indicates that the explosive loading was sufficient to obtain a wavy interface. The wave formation in explosive welding can be regarded as a special case of general phenomena of interfacial wave formation under certain flow circumstance. As reported by Wronka, the collision area was subjected to an axial force normal to plate surface and a tangent force parallel to plate surface during explosive welding. Given that the collision pressure p was greater than the theoretical yield point IJmax and the slowest velocity of forming waves Vf less than the sound velocity in the metal, the effective wavy joints would be formed [29]. Wave formation appears to be the result of variations in the velocity distribution at collision point and periodic disturbances of materials [30]. In this study, the wavy length and the amplitude of 2205/X65 interface were approximately 1.3 mm and 0.6 mm, respectively. Fig. 2 shows the microstructure of the cross-section of the bimetallic sheet at 7 typical positions. The 7 typical positions are indicated by the rectangles B, C, D, E, F 6

and G in Fig. 2a and the rectangle H in Fig. 2d, respectively, and the corresponding microstructure is shown in Fig. 2b, c, d, e, f, g and h. Fig. 2a shows that the welded interface could be identified by a characteristic sharp transition between two materials, but local melted zones were also encountered in the front or back slope of waves in the interface. According to Fig. 2a, two types of bond were encountered at the 2205/X65 interface in this study, namely metal±metal and metal±solidified melt bond. This is consistent with the results of Kacar and co-workers. [27]. Formation of local melted zones was attributed to the adiabatic heating of trapped jet inside vortices at the front or back slope of waves as a result of density difference or the adiabatic heating of gases compressed between the plates [30]. In the localized melted region with a small width, the small amount of liquid metal and the rapid cooling speed resulted in a fine equiaxed grain structure, as shown in Fig. 2b, d and h. From Fig. 2b and d, it can be seen that in the localized melted region with a large width, a coarse columnar crystal structure formed along the perpendicular direction to the interface. At the final solidification site where the columnar crystal grains met, shrinkage and porosity were observed frequently. In Kacar et al.¶s study, fine equiaxed grain structure was also observed in the local melted zone between 2205 stainless steel and X65 steel, but coarse columnar crystal structure was absent in the local melted zone [27]. This can be explained by the fact that the maximum width of local melted zone obtained in Kacar R et al.¶s study and our work was about 10 ȝPDQG100 ȝP, respectively. Fig. 2b also shows that the metal from the X65 side extended into the 2205 stainless steel, forming a peninsula-like morphology, as indicated by the arrow in Fig. 2b. The peninsula-like morphology resulted from the combined action of the detonation force and the metal vortex flow. When the detonation force and the metal vortex flow were very intense, an island-like morphology was also likely to form: the arrow shown in Fig. 2a shows this morphology. Fig. 2b and d also show that the fibrous morphology along the interface direction was significantly elongated at the stainless steel side near the interface. This result was consistent with the previous studies [19-26]. Fig. 2c and e show the microstructure near the front and back slopes of the wavy interface, with a very clear demarcation between the two materials. Dramatic non-uniform deformation occurred in the 2205 metal near the interface, whereas the initial microstructure of the stainless steel side completely disappeared. The huge pressure and shear force of the explosion welding process on the duplex stainless steel side near the interface produced many thin strip morphology that extended out from the interface, which were commonly known as fly line structures or shear bands. Fig. 2f shows that a dramatic plastic deformation occurred in the austenitic stainless steel above the fly line, which appeared as a flocculent morphology. A small degree of plastic deformation in the region below the fly line better preserved the original 7

microstructure of 2205 stainless steel. Fig. 2g shows that the original microstructure was essentially retained in the stainless steel far away from the interface. The aforementioned analysis reveals a highly uneven structure near the interface that resulted from the severe deformation of the interface and the presence of the morphologies of the localized melted region, the peninsula, and the island. In addition, solidification cracks, pores and other defects were observed near the interface. These phenomena would significantly influence the subsequent pipe manufacturing processes of the bimetallic sheet, such as bending, plate rolling and tube diameter expanding. 3.2 SEM observation of the interface Fig. 3a shows four typical positions (i.e., the rectangular areas b, c, d, and e in Fig. 3a) around the wavy interface. The SEM observation results of the morphologies around these four positions are shown in Fig. 3b, c, d, and e. Fine crystal grains ranging from 0.5-2 microns in size were observed all over the narrow localized melted region around the peak, as shown in Fig. 3b and d. Fig. 3c and e show a clear demarcation line between the two materials near the rising and falling edges of the wavy interface. The austenite phase in the 2205 region adjacent to the interface was elongated, and a large quantity of crystal grains was distributed within the austenite region, as shown in Fig. 3e. An elongated fibrous structure was observed in the X65 region farther from the interface. No elongated fibrous structure was observed in the X65 region adjacent to the interface, most likely because recovery and recrystallization occurred in this region, as shown in Fig. 2b and e. Similar result was reported by Kacar R et al., who studied the microstructure and micro-hardness of 2205 steel/steel bimetallic sheet. They observed a band between the local melted zone and the flyer plate and suggested that formation of such a band was related to the heat generation in the interface region that caused annealing of the shock-hardened material in 2205 side. The authors noted that Kacar et al. measured the maximum mcro-hardness about 25 ȝP away from interface in both flyer plate and base plate, which indicated that annealed band was formed not only in the flyer plate but also in the base plate [27]. 3.3 Distribution of alloy elements across interface Fig. 4a shows five positions near the interface for the electron probe linear scanning analysis (i.e., lines b, c, d, e, and f in Fig. 4a). Fig. 4f shows a high Cr content in the metal inside the melting region for the 2205 side near the localized melted region with a large width, whereas the X65 side had a lower Cr content. The width of the region over which the Cr content changed significantly was about 2ȝP. Fig. 4c and e show that the region over which there was a significant change in Cr content near the rising and falling edges of the wavy interface was approximately 2-3 microns wide. Overall, the diffusion layer of the alloying elements near the interface of the bimetallic sheet was relatively thin; however, the distribution of the 8

two materials was staggered in the localized melted region near the peak and trough of wavy interface, resulting in a more complex distribution of the alloying elements. Fig. 4b and d show the significant fluctuations in Cr and Fe contents near the localized melted region with a small width. Vigorous detonation forces and a metal vortex flow could have caused a portion of the liquid or semi-solid X65 metal to enter the 2205 side and be rapidly cooled and solidified, producing the results shown in Fig. 4b and d. Figs. 5 and 6 show the detailed test results for the alloy element composition near the crest and trough. 3.4 Mechanical properties 3.4.1 Micro-hardness Fig. 7a and b show the test position and results of the Vickers hardness test near the interface, from which a remarkable difference in the micro-hardness between 2205 and X65 can be observed. Fig. 7b demonstrates that the micro-hardness generally decreased as the distance from the welding interface increased. Similar results have been reported by many researchers [31]. It is generally accepted that this could be attributed to the high plastic deformation in the welding zone relative to the area further away. Fig. 7b also shows that the micro-hardness did not change substantially in either metal if the distance from the interface was greater than 0.5 mm. The maximum Vickers hardness of the X65 near the interface was approximately 300 HV, which was approximately 33% higher than the 225 HV hardness of the X65 far away from the interfacial region. The maximum Vickers hardness of the 2205 near the interface was approximately 470 HV, which was approximately 24% higher than the hardness of the 2205 of 380 HV far away from the interfacial region. It is also seen from Fig. 7b that the maximum micro-hardness of 2205 material on both line 1 and line 2 was measured about 0.2mm away from 2205/X65 interface, which was inconsistent with the frequently reported result that the maximum micro-hardness was detected near the interface. This phenomenon might be explained by the fact that the microstructure in the 2205 region adjacent to the interface was quite different from that in the region far away from interface, as shown in Fig. 3e and Fig. 4e. Kacar et al. who studied explosively welded 2205/vessel steel sheets observed similar phenomenon and suggested that this might due to the heat generation in the interface region that caused annealing of the shock-hardened material [27]. Noting that annealed band in austenitic stainless steel/carbon steel bimetallic sheet was rarely reported, the authors would like to add that this phenomenon may also related to the higher yield strength of 2205 stainless steel compared to standard austenitic stainless steel, which means that more explosive should be used in the welding process and therefore would lead to higher heat generation. 3.4.2 Stratified tensile test

9

The micro-hardness test results in Fig. 7b show that the mechanical properties of the bimetallic sheet dramatically changed in the thickness direction; thus, a stratified tensile test was carried out for the bimetallic sheet. The stratified specimens were collected using the method shown in Fig. 8, and the tensile specimens were then fabricated to the dimensions shown in Fig. 9, where the stretching direction was consistent with the direction of the detonation wave. A tensile test was also carried out for the 2205 and X65 plates that did not undergo explosive welding. Fig. 10 shows the stress-strain curve obtained from the tensile test. Without explosion welding, the stress-strain curves of the 2205 base metal and the X65 base metal obtained from the tensile test were in good agreement with the literatures [32, 33]. Fig. 11 summarizes the pattern of variation in the mechanical properties in the thickness direction of the bimetallic sheet and the variation in microstructure along the thickness direction. The results showed that the tensile strength of the material in the second layer was slightly lower than that of the material in the first layer and that the plasticity of the material in the second layer was relatively poor. The test results showed that after the explosion welding, the tensile strength of the first layer (i.e., stainless steel) increased from about 755 MPa to 1137 MPa, the tensile strength of the second layer (i.e., near the interface) was approximately 1080 MPa, the tensile strength of the third layer of X65 increased from 568 MPa to 652 MPa. The results indicate that the strength of second layer was higher than the average strength between the first layer and the second layer. This was due to that the second layer suffered the most severe deformation in explosive welding, as shown in Fig. 11b, c and d. According to Fig. 11d, e, f, g and h, plastic deformation in base plate decreased with the increase of the distance to 2205/X65 interface. Therefore, with the increase of the distance to 2205/X65 interface, tensile strength of stratified material in X65 side decreased and elongation rate of stratified material in X65 side increased. The mechanical properties from layers 4 to 6 were essentially unchanged at a tensile strength of approximately 630 MPa. Hashemi et al. showed that the relationship between the tensile strength (UTS) and the hardness (HV) of X65 pipe steel could be described by the following equation [34]:

UTS 1.3u HV  344

(1).

In the present study, the hardness of X65 in layers 4 to 6 away from the interface was approximately 225 HV. The tensile strength calculated using the equation above was about 636 MPa. The tensile strength of the material in layer 6 was measured using the tensile test at 630 MPa, which was in good agreement with the result of Hashemi and co-workers. 3.4.3

Tensile shear test

A tensile shear test was conducted for the 2205/X65 bimetallic sheet following GB/T6396-2008 [28]. The specimen dimensions are shown in Fig. 12. Fig. 13 shows the three specimens before and after the tensile shear test was conducted. Fig. 14 10

shows the force-distance curve obtained from the tensile shear test and the relative position of the corresponding fracture. The average shear fracture strength for the three specimens was about 470 MPa, which complied with the GB/T6396-2008 standard [28]. 3.4.4

Bending test

The 2205/X65 bimetallic sheet prepared in this study will be used to manufacture oil and gas pipelines, where the inner and outer pipes should be 2205 stainless steel and X65 pipe steel, respectively. Therefore, a three-point face bending test was applied to the explosively welded 2205/X65 bimetallic sheet, and the bended specimens are shown in Fig. 15. The total thickness of the bending specimen was 8 mm, wherein the thickness of the stainless steel layer was 2 mm, and the thickness of the X65 layer was 6 mm. In the face bending test, the indenter diameter was 40 mm, the cylinder diameter was 30 mm, and the distance between the axes of the two cylinders was 94 mm. Fig. 15 shows that after bending up to 180 degrees, the specimens were neither separated nor cracked, indicating a reliable joining between 2205 and X65. JCOE pipe is manufactured by repeatedly bending a flat plate to form a tube. To determine the weak position near the interface of the 2205/X65 bimetallic sheet in the bending process, 45° face bending tests and 45° root bending tests were conducted on 18-mm thick specimens using an indenter with a small diameter (i.e., 25 mm) and a roller span of 120 mm, as shown in Fig. 16a. Fig. 16b shows the specimens after bending. The longitudinal section (i.e., parallel to the lengthwise direction of the specimen) at the portion of the specimen with the maximum curvature was ground, polished and etched before observation under an optical microscope. Fig. 16c and d are the respective optical micrographs of the longitudinal sections for the 45° root bending and 45°face bending specimens near the interface, which show that cracks did not occur in the interface after the bending test, whereas voids appeared in the peninsula and island morphologies near the interface. This can be partly explained by the results in Fig. 2, which demonstrates that shrinkage, crack and porosity were frequently observed near the peninsula and island morphologies. Fig. 16c also shows numerous dispersed micropores in the X65 side near the interface of a root bending specimen. 3.5

SEM fracture analysis

3.5.1 Fracture in stratified tensile specimen Fig. 17a shows the macroscopic fracture morphology of the first layer of material (i.e., 2205 stainless steel). Fig. 17b, e, and f show high resolution fractography of the positions B, E and F in Fig.17a, respectively. Fig. 17b shows numerous small dimples and a few large dimples in position B, and Fig. 17d shows the high resolution morphology of one of these large dimples. Fig. 17d shows that the 11

large dimples were oval, and the major axis direction is shown by the dotted line in the image. Fig. 17b also shows multiple layered cracks (see the arrow in Fig. 17b). The crack directions of these layered cracks were in the direction of the long axes of the large dimples. It seems that the large oval dimples and the layered crack formation might relate to the elongated fibrous morphology in the 2205 layer (see the top half of Fig. 17c). Fig. 17e and f show that the dimples were distributed throughout positions E and F in the macroscopic fracture, and shearing dimple morphology was observed at position F near the edge of the macroscopic fracture. Observation of the optical structure of the cross-section of the bimetallic sheet showed that a columnar grain structure formed in the localized melted zone with a large width, and cracks and pores easily formed at the intersections of the columnar grains, as shown in Fig. 18a and Fig. 2a. The SEM analysis of the tensile fracture revealed the quasi-cleavage morphology at the fracture of the columnar grain structure in the localized melted zone, as shown in Fig. 18b. Fig. 18c shows high resolution microscopic fractography of position C in Fig. 18b. Obviously, the presence of a localized melted zone could weaken the mechanical properties of the bimetallic sheet, as reported by Kacar R et al. [19] and Akbari Mousavi SAA et al. [23]. Fig. 18d, e, f, and g show the target sites of EDS analysis and the corresponding results. The upper part of Fig. 18d (i.e., rectangle 3 in Fig. 18d) consists of the X65 material, and the lower part (i.e., rectangles 5 and 6 in Fig. 18d) shows the 2205 material. Fig. 19 also shows the fractography of the tensile fracture in the localized melted zone of the second layer, in which Fig. 19b and c show high resolution fractography of positions B and C in Fig. 19a. It can be seen from Fig. 19 that a quasi-cleavage fracture surface formed in the columnar crystal zone near the interface. Fig. 20a shows a Y-shaped crack in the tensile fracture of layer 2. Fig. 20b shows high resolution fractography of position B in Fig. 20a. Fig. 20a shows three regions with distinct microscopic morphologies. The upper area exhibited dimple morphology, the lower area exhibited quasi-cleavage fracture morphology, and small dimples were distributed throughout the area in between these two regions (see Fig. 20b). The EDS detection results (Fig. 20d, e, f, and g) showed that the upper area in Fig. 20a was in the 2205 region, whereas the middle and lower areas were in the X65 region. Fig. 21 shows another example of a Y-shaped crack in the tensile fracture of layer 2. The Y-shaped cracks in Fig. 20 and 21 may related to the concentrated stress that easily occurred at the intersection of three different structures (see Fig. 20c). Fig. 22 shows the fracture morphology of layer 6 (i.e., for X65 far away from the interface), which appears as a dimple fracture. Fig. 22a shows the macroscopic fracture morphology, and Fig. 22b, c and d shows the microscopic morphologies of the positions B, C and D in Fig. 22a, respectively. Fig. 22b shows a dimple microscopic morphology at position B of the macroscopic fracture, and some large dimples were distributed among numerous small dimples. Spiral corrugation was observed on the inner surface of the large dimples. Fig. 22c shows that a dimple-like 12

microscopic morphology was also observed near position C of the macroscopic fracture, and small dimples were distributed throughout this area. Fig. 22d shows a shearing dimple microscopic morphology at position D near the edge of the macroscopic fracture. The fracture morphology of the material in the third layer (i.e., X65 near the interface) was very similar to that of the material in the sixth layer, which both corresponded to dimple-type fractures. The difference between these morphologies was that there were fewer large dimples in the center of the fracture in the third layer than the sixth layer, indicating that the plasticity of the material in the third layer was worse than that of the sixth layer. Fig. 23a and b show the microscopic morphology near the center of the macroscopic fracture for the materials in the third layer and the sixth layer. 3.5.2 SEM analysis of tensile shear fracture The tensile shear tests demonstrated that the explosion-bonded 2205/X65 bimetallic sheet had good shear strength and that shearing fracture primarily occurred for the dimple morphology. Fig. 24a shows the macro-morphology of the tensile shear fracture in the 2205/X65 bimetallic sheet, Fig. 24b shows the distribution of alloying elements near position B of the macro fracture, and Fig. 24c shows the high resolution microscopic morphology near position B. The distribution of alloying elements in Fig. 24b and e showed that position B was near the interface between 2205 and X65 regions. Fig. 24d shows high resolution fractography of position D in Fig. 24b (i.e., the inner wall of the pores). The EDS analysis revealed that the Cr content in the inner wall of the pores was approximately 8.91%, indicating that the metal around the pores was a mixture of the two materials, which implied that the pores were located in the localized melted zone. The EDS analysis showed that most of the tensile shear fracture occurred in the X65 material. Only a very small part of the fracture occurred in the 2205 region, which was accompanied by defects such as pores. It seems that fracture path in tensile shear test was dependent on the strength difference between flyer plate and base plate. In this work, the strength of flyer plate was much higher than that of base plate, as shown in Fig. 10. Acarer et al. reported that fracture was obtained at either upper or lower plate of the bimetallic sheet during tensile shear test other than at the explosively joined area [25]. In Acarer et al.¶s study, the strength of flyer plate was similar to that of the base plate. In summary, the interface of explosively welded 2205/X65 bimetallic sheet exhibited reliable and robust shear strength and bending properties. However, severe inhomogeneity was observed in the microstructure and mechanical properties of the 2205/X65 bimetallic sheet, and defects such as pores, localized melted zone, peninsula morphology and island morphology were observed near the interface. Impacts of these characteristics on the subsequent JCOE forming process of 2205/X65 bimetallic pipe should be studied. 13

4. Conclusions In this work, the microstructure and mechanical properties of explosive welded 2205 stainless steel/X65 pipe steel bimetallic sheet which would be used to manufacture JCOE welded oil and gas pipe were investigated, with emphasis on the inhomogeneity of both microstructure and mechanical properties of the bimetallic sheet obtained. The following conclusions can be drawn from this study: (1) Large scale 2205/X65 bimetallic sheets were obtained by explosive welding. Two types of bond were observed at wavy 2205/X65 interface, namely metal±metal and metal±solidified melt. Numerous fine grains ranging in size from 0.5-2 microns were observed in a localized melted zone with a small width; coarse columnar crystal structures growing along the perpendicular direction to the interface were observed in the localized melted region with a large width. Shrinkage, porosity and other defects easily occurred at the intersecting area of the columnar crystals. (2) The X65 region far away from the interface has a significantly elongated structure, whereas a fibrous structure was not observed in the X65 region adjacent to the interface most likely because recovery and recrystallization occurred in this region. In the 2205 region adjacent to the interface, numerous fine grains were distributed over elongated austenite phase. (3) The fractures of the stratified tensile specimens in the stainless steel layer and the X65 layer of the bimetallic sheet exhibited dimple morphology, and multiple layered cracks were observed in the fracture in the stainless steel layer of the material. The tensile fracture near the 2205/X65 interface primarily occurred in the dimple morphology, with few quasi-cleavage fracture areas that were prone to occur in the localized melted zone containing coarse columnar crystals. (4) A stratified tensile test was conducted to determine the variation in the mechanical properties along the thickness direction of the bimetallic sheet. (5) The tensile shear strength of the 2205/X65 bimetallic sheet met the requirements of ASME S264. Tensile shear fracture primarily had dimple morphology. Tensile shear fracture primarily occurred inside the X65 material. A very small part of the fracture occurred in the 2205 region and was accompanied by defects such as pores. (6) Under tensile load, a typical Y-shaped crack was prone to occur in the junction of three different structural regions (i.e., the elongated structure in 2205, the elongated structure in X65, and the recrystallized structure region or the local melted region). (7) Under bending load, void was prone to appeared in the peninsula and island morphologies near the interface. Acknowledgements This work was supported by the National High Technology Research and Development Program of China (Grant No. 2013AA031303HZ). 14

References [1] Miao GH, Ma HH, Shen ZW, Yu Y. Research on honeycomb structure explosives and double sided explosive cladding. Mater Des 2014; 63: 538±43. [2] Guo X, Tao J, Wang W, Li H, Wang C. Effects of the inner mould material on the aluminium±316L stainless steel explosive clad pipe. Mater Des 2013; 49: 116±22. [3] Zhou N, Wang JX, Peng CC, Tong ZB. Experimental and numerical study on the ballistic resistance of steel±fibre reinforced two±layer explosively welded plates. Mater Des 2014; 54: 104±11. [4] Ahmet D, Gülenc B, Findik F. Examination of copper/stainless steel joints formed by explosive welding. Mater Des 2005; 26: 497±507. [5] Bina MH, Dehghani F, Salimi M. Effect of heat treatment on bonding interface in explosive welded copper/stainless steel. Mater Des 2013; 45: 504±9. [6] Raghukandan K. Analysis of the explosive cladding of Cu±low carbon steel plates. J Mater Process Technol 2003; 139: 573±7. [7] Gülenc B. Investigation of interface properties and weldability of aluminum and copper plates by explosive welding method. Mater Des 2008; 29: 275±8. [8] Kahraman N, Gulenc B. Microstructural and mechanical properties of Cu±Ti plates bonded through explosive welding process. J Mater Process Technol 2005; 169: 67±71. [9] Xia H, Wang S, Ben H. Microstructure and mechanical properties of Ti/Al explosive cladding. Mater Des 2014; 56: 1014±19. [10] Bataev IA, Bataev AA, Mali VI, Pavliukova DV. Structural and mechanical properties of metallic±intermetallic laminate composites produced by explosive welding and annealing. Mater Des 2012; 35: 225±34. [11] Nizamettin K, Gülenc B, Findik F. Joining of titanium/stainless steel by explosive welding and effect on interface. J Mater Process Technol 2005; 169: 127±33. [12] Song J, Kostka A, Veehmayer M, Raabe D. Hierarchical microstructure of explosive joints: example of titanium to steel cladding. Mater Sci Eng: A 2011; 528: 2641±7.

15

[13] Wang B, Chen W, Li J, Liu Z, Zhu X. Microstructure and formation of melting zone in the interface of Ti/NiCr explosive cladding bar. Mater Des 2013; 47: 74±79. [14] Manikandan P, Hokamoto K, Fujita M, Raghukandan K, Tomoshige R. Control of energetic conditions by employing interlayer of different thickness for explosive welding of titanium/304 stainless steel. J Mater Process Technol 2008; 195: 232±40. [15] Liu W, Liu K, Chen Q, Wang J, Yan H, Li X. Metallic glass coating on metals plate by adjusted explosive welding technique. Appl Surf Sci 2009; 255: 9343±7. [16] Zareie Rajani HR, Akbari Mousavi SAA, Madani Sani F. Comparison of corrosion behavior between fusion cladded and explosive cladded Inconel 625/plain carbon steel bimetal plates. Mater Des 2013; 43: 467±74. [17] Gerland M, Presles HN, Guin JP, Bertheau D. Explosive cladding of a thin Nifilm to an aluminium alloy. Mater Sci Eng 2000; A280: 311. [18] Gong S, Li Z, Xiao Z, Zheng F. Microstructure and property of the composite laminate cladded by explosive welding of CuAlMn shape memory alloy and QBe2 alloy. Mater Des 2009; 30: 1404±8 [19] Kacar R, Acarer M. An investigation on the explosive cladding of 316L stainless steel±din p355gh steel. J Mater Process Technol 2004; 152: 91±6. [20] Kaya Y, Nizamettin K. An investigation into the explosive welding/cladding of Grade A ship steel/AISI 316L austenitic stainless steel. Mater Des 2013; 52: 367±72. [21] Mendes R, Ribeiro JB, Loureiro A. Effect of explosive characteristics on the explosive welding of stainless steel to carbon steel in cylindrical configuration. Mater Des 2013; 51: 182±92. [22] Zamani E, Liaghat GH. Explosive welding of stainless steel±carbon steel coaxial pipes. J Mater Sci 2012; 47: 685±95. [23] Akbari±Mousavi SAA, Al±Hassani STS, Atkins AG. Bond strength of explosively welded specimens. Mater Des 2008; 29: 1334±52. [24] Akbari±Mousavi SAA, Al-Hassani STS. Finite element simulation of explosively±driven plate impact with application to explosive welding. Mater Des 2008; 29: 1±19.

16

[25] Acarer M, Gulenc B, Findik F. The influence of some factors on steel/steel bonding quality on their characteristics of explosive welding joints. J Mater Sci 2004; 39: 6457±66. [26] Acarer M, Gülenc B, Findik F. Investigation of explosive welding parameters and their effects on microhardness and shear strength. Mater Des 2003; 24: 659±64. [27] Kacar R, Acarer M. Microstructure±property relationship in explosively welded duplex stainless steel±steel. Mater Sci Eng A 2003; 363: 290±6. [28] GB/T6396±2008. Clad steel plates±Mechanical and technological test, China National Standardization Management Committee; May 13, 2008. [29] Wronka B. Testing of explosive welding and welded joints: joint mechanism and properties of explosive welded joints. J Mater Sci 2010; 45: 4078±83. [30] Mousawi SAAA, Sartangi PF. Experimental investigation of explosive welding of cp±titanium/AISI 304 stainless steel. Mater Des 2009; 30: 459±68. [31] Findik F. Recent Developments in Explosive Welding. Mater Des 2011; 32: 1081±93. [32] Tsai WT, Chen MS. Stress corrosion cracking behavior of 2205 duplex stainless steel in concentrated NaCl solution. Corros Sci 2000; 42: 545±59. [33] Fukuda N, Hagiwara N, Masuda T. Effect of prestrain on tensile and fracture toughness properties of line pipes. J Offshore Mech Arct 2005; 127: 263±8. [34] Hashemi SH. Strength±hardness statistical correlation in API X65 steel. Mater Sci Eng: A 2011; 528: 1648±55.

17

List of Table Captions

Table 1 Chemical composition of X65 pipe steel and 2205 duplex stainless steel (wt %)

18

Table 1 Chemical composition of X65 pipe steel and 2205 duplex stainless steel (wt %) C

Si

Cr

Mn

Ni

Cu

Mo

V

P

S

Fe

X65

0.053

0.33

0.07

1.18

0.16

0.14

0.06

0.031

0.016

İ0.005

Balance

2205

0.021

0.56

22.59

1.13

5.29

-

3.45

-

0.018

0.001

Balance

19

List of Figure Captions

Fig. 1 Cross-section of 2205/X65 bimetallic sheet achieved by explosive welding Fig. 2 Optical microscopy images of interface in explosion-bonded 2205/X65 bimetallic sheet (Fig. 2h is a SEM image): (a) wavy interface, (b) high resolution image (HRI) of position B in Fig. 2a, (c) HRI of position C in Fig. 2a, (d) HRI of position D in Fig. 2a, (e) HRI of position E in Fig. 2a, (f) HRI of position F in Fig. 2a, (g) HRI of position G in Fig. 2a, and (h) HRI of position H in Fig. 2d Fig. 3 SEM images of the interface in an explosively welded 2205/X65 bimetallic sheet: (a) wavy interface, (b) high resolution image (HRI) of position B in Fig. 3a, (c) HRI of position C in Fig. 3a, (d) HRI of position D in Fig. 3a, and (e) HRI of position E in Fig.3a Fig. 4 Distribution of elements across the 2205/X65 interface: (a) wavy interface, (b) distribution of alloy elements (DAE) along line b in Fig. 4a, (c) DAE along line c in Fig. 4a, (d) DAE along line d in Fig. 4a, (e) DAE along line e in Fig. 4a, and (f) DAE along line f in Fig. 4a Fig. 5 EDS analysis results of local melted zone at 2205 wave crest: (a) results for rectangle 1 in Fig. 5e, (b) results for rectangle 2 in Fig. 5e, (c) results for rectangle 3 in Fig. 5e, (d) results for rectangle 4 in Fig. 5e, and (e) target areas of EDS analysis Fig. 6 EDS analysis results of local melted zone at 2205 wave trough: (a) results for rectangle 1 in Fig. 6e, (b) results for rectangle 2 in Fig. 6e, (c) results for rectangle 3 in Fig. 6e, (d) results for rectangle 4 in Fig. 6e, and (e) target areas of EDS analysis Fig. 7 Micro-hardness profile across interface Fig. 8 Stratified scheme for stratified tensile test for bimetallic sheet Fig. 9 Dimensions of stratified tensile specimen (Unit: mm) Fig. 10 Stress-strain curves of stratified tensile test for bimetallic sheet and tensile test for base metal Fig. 11 Changes in mechanical properties and microstructures along the thickness direction of 2205/X65 bimetallic sheet Fig. 12 Schematic showing specimen dimensions for tensile shear test following GB/T6396-2008 standard (units: mm) [28] Fig. 13 Specimens (a) before tensile shear test and (b) after tensile shear test 20

Fig. 14 Force-distance curve and position of tensile shear fracture relative to interface Fig. 15 Three-point face bending test: (a) face bending test and (b) bended specimens Fig. 16 Bending test under extreme conditions to observe weakest point: (a) three-point bending test, (b) bended specimens, (c) voids in root-bended specimen, and (d) voids in face-bended specimen Fig. 17 Fractography of Layer 1 (2205) after tensile test: (a) macro-fractography of Layer 1, (b) high resolution fractography (HRF) of position B in Fig. 17a, (c) Schematic of elongated optical structure in 2205 side, (d) HRF of position D in Fig. 17b, (e) HRF of position E in Fig. 17a, and (f) HRF of position F in Fig. 17a Fig. 18 Fractography of the local melted-solidified zone consisting of columnar grains and porosity: (a) optical image of local melted zone showing porosity, (b) fractography of local melted zone showing porosity, (c) high resolution fractography of position C in Fig. 18b, (d) target areas of EDS analysis, (e) EDS analysis result (EAR) for rectangle 3 in Fig. 18d, (f) EAR for rectangle 5 in Fig. 18d, and (g) EAR for rectangle 6 in Fig.18d Fig. 19 Fractography of localized melted/solidified zone consisting of columnar grains: (a) fractography of local melted/solidified zone consisting of columnar grains; (b) high resolution fractography (HRF) of position B in Fig. 19a; and (c) HRF of position C in Fig. 19a Fig. 20 Y-shaped crack at the intersection of three regions with different structures: (a) SEM image of Y-shaped crack; (b) high resolution fractography of position B in Fig. 20a; (c) optical image of intersection of three regions with different structures; (d) target areas of EDS analysis; (e) EDS analysis results (EAR) for rectangle 1 in Fig. 20d; (f) EAR for rectangle 2 in Fig. 20d; and (g) EAR for rectangle 3 in Fig. 20d Fig. 21 Example of Y-shaped crack in fracture of layer 2 Fig. 22 Fractography of layer 6 (X65 far away from interface): (a) macro-fractography; (b) high resolution fractography (HRF) of position B in Fig. 22a; (c) HRF of position C in Fig. 22a; and (d) HRF of position D in Fig. 22a Fig. 23 Comparison of fractography between (a) layer 3, X65 near to interface and (b) layer 6, X65 far away from interface Fig. 24 Fractography of tensile shear test and EDS analysis of fracture: (a) macro-fracture; (b) distribution of alloy elements in position B in Fig. 24a and target areas of EDS analysis;

21

(c) high resolution fractography of position C in Fig. 24b; (d) high resolution fractography of position D in Fig. 24b; and (e) EDS analysis results

22

Fig. 1 Cross-section of 2205/X65 bimetallic sheet achieved by explosive welding

23

Fig. 2 Optical microscopy images of interface in explosion-bonded 2205/X65 bimetallic sheet (Fig. 2h is a SEM image): (a) wavy interface, (b) high resolution image (HRI) of position B in Fig. 2a, (c) HRI of position C in Fig. 2a, (d) HRI of position D in Fig. 2a, (e) HRI of position E in Fig. 2a, (f) HRI of position F in Fig. 2a, (g) HRI of position G in Fig. 2a, and (h) HRI of position H in Fig. 2d

24

Fig. 3 SEM images of the interface in an explosively welded 2205/X65 bimetallic sheet: (a) wavy interface, (b) high resolution image (HRI) of position B in Fig. 3a, (c) HRI of position C in Fig. 3a, (d) HRI of position D in Fig. 3a, and (e) HRI of position E in Fig. 3a

25

Fig. 4 Distribution of elements across the 2205/X65 interface: (a) wavy interface, (b) distribution of alloy elements (DAE) along line b in Fig. 4a, (c) DAE along line c in Fig. 4a, (d) DAE along line d in Fig. 4a, (e) DAE along line e in Fig. 4a, and (f) DAE along line f in Fig. 4a

26

Fig. 5 EDS analysis results of local melted zone at 2205 wave crest: (a) results for rectangle 1 in Fig. 5e, (b) results for rectangle 2 in Fig. 5e, (c) results for rectangle 3 in Fig. 5e, (d) results for rectangle 4 in Fig. 5e, and (e) target areas of EDS analysis

27

Fig. 6 EDS analysis results of local melted zone at 2205 wave trough: (a) results for rectangle 1 in Fig. 6e, (b) results for rectangle 2 in Fig. 6e, (c) results for rectangle 3 in Fig. 6e, (d) results for rectangle 4 in Fig. 6e, and (e) target areas of EDS analysis

28

Fig. 7 Micro-hardness profile across interface

29

Fig. 8 Stratified scheme for stratified tensile test for bimetallic sheet

30

Fig. 9 Dimensions of stratified tensile specimen (Unit: mm)

31

Fig. 10 Stress-strain curves of stratified tensile test for bimetallic sheet and tensile test for base metal

32

Fig. 11 Changes in mechanical properties and microstructures along the thickness direction of 2205/X65 bimetallic sheet.

33

Fig. 12 Schematic showing specimen dimensions for tensile shear test following GB/T6396-2008 standard (units: mm) [28]

34

Fig. 13 Specimens (a) before tensile shear test and (b) after tensile shear test

35

Fig. 14 Force-distance curve and position of tensile shear fracture relative to interface

36

Fig. 15 Three-point face bending test: (a) face bending test and (b) bended specimens

37

Fig. 16 Bending test under extreme conditions to observe weakest point: (a) three-point bending test, (b) bended specimens, (c) voids in root-bended specimen, and (d) voids in face-bended specimen

38

Fig. 17 Fractography of Layer 1 (2205) after tensile test: (a) macro-fractography of Layer 1, (b) high resolution fractography (HRF) of position B in Fig. 17a, (c) Schematic of elongated optical structure in 2205 side, (d) HRF of position D in Fig. 17b, (e) HRF of position E in Fig. 17a, and (f) HRF of position F in Fig. 17a

39

Fig. 18 Fractography of the local melted-solidified zone consisting of columnar grains and porosity: (a) optical image of local melted zone showing porosity, (b) fractography of local melted zone showing porosity, (c) high resolution fractography of position C in Fig. 18b, (d) target areas of EDS analysis, (e) EDS analysis result (EAR) for rectangle 3 in Fig. 18d, (f) EAR for rectangle 5 in Fig. 18d, and (g) EAR for rectangle 6 in Fig. 18d

40

Fig. 19 Fractography of localized melted/solidified zone consisting of columnar grains: (a) fractography of local melted/solidified zone consisting of columnar grains; (b) high resolution fractography (HRF) of position B in Fig. 19a; and (c) HRF of position C in Fig. 19a

41

Fig. 20 Y-shaped crack at the intersection of three regions with different structures: (a) SEM image of Y-shaped crack; (b) high resolution fractography of position B in Fig. 20a; (c) optical image of intersection of three regions with different structures; (d) target areas of EDS analysis; (e) EDS analysis results (EAR) for rectangle 1 in Fig. 20d; (f) EAR for rectangle 2 in Fig. 20d; and (g) EAR for rectangle 3 in Fig. 20d

42

Fig. 21 Example of Y-shaped crack in fracture of layer 2

43

Fig. 22 Fractography of layer 6 (X65 far away from interface): (a) macro-fractography; (b) high resolution fractography (HRF) of position B in Fig. 22a; (c) HRF of position C in Fig. 22a; and (d) HRF of position D in Fig. 22a

44

Fig. 23 Comparison of the fractography between: (a) layer 3, X65 near to interface and (b) layer 6, X65 far away from interface

45

Fig. 24 Fractography of tensile shear test and EDS analysis of fracture: (a) macro-fracture; (b) distribution of alloy elements in position B in Fig. 24a and target areas of EDS analysis; (c) high resolution fractography of position C in Fig. 24b; (d) high resolution fractography of position D in Fig. 24b; and (e) EDS analysis results

46

Highlights 2205/X65 bimetallic plate was produced

by explosive welding.

Microstructural inhomogeneity of 2205/X65 bimetallic sheet was investigated. Mechanical inhomogeneity of 2205/X65 bimetallic sheet was studied. Distribution of alloy elements across 2205/X65 interface was studied. Weak positions of 2205/X65 sheet in bending process were revealed.

47